Seismic analyses of conventional and improved marginal wharves

EARTHQUAKE ENGINEERING AND STRUCTURAL DYNAMICSEarthquake Engng Struct. Dyn. 2011; 00:1–19Published online in Wiley InterScience (www.interscience.wiley.com). DOI: 10.1002/eqe

Seismic Analyses of Conventional and Improved MarginalWharves

Maurizio M. Chiaramonte∗†, Pedro Arduino‡, Dawn E. Lehman§, Charles W. Roeder‡

Department of Civil and Environmental Engineering, University of Washington, Seattle, WA 98195, U.S.A.

SUMMARY

Marginal wharves are key components in providing functionality of port facilities. Ports are centralcomponents of the United States economy. Earthquake damage to a port can disrupt the economic stability.Therefore port facilities must be able to quickly return to full operation shortly after a seismic event. Priorstudies have shown that integrity of marginal wharves may be compromised by excessive soil movementand structural damage. The latter is often localized at pile-to-wharf connections and in the pile body buriedwithin the soil. Recent research has resulted in an improved connection design which mitigates damage.This study was undertaken to evaluate the full seismic performance of marginal wharves including bothconventional and damage resisting connections. A series of finite element (FE) models of a representativepile supported wharf facility were created. The models varied in their moment resisting pile-to-wharfconnections. A total stress analysis approach was used to capture the soil response along with p-y, t-zand Q-z soil-structure interaction (SSI) springs. Validated connection interface elements were integratedwith non-linear frame elements to simulate the marginal wharf structure and substructure. Non-linear staticpushover and dynamic time history analyses, for three different hazard levels, were performed. The resultsof the numerical simulations were used to assess the performance of the marginal wharf including estimatesof crane damage and port downtime. Copyright c© 2011 John Wiley & Sons, Ltd.

Received . . .

KEY WORDS: Marginal Wharves; Ports; Soil-Structure Interaction; Performance Based Design;Earthquake Engineering.

1. INTRODUCTION

Seaports are vital components of economic stability. In the United States alone portfacilities are responsible for the timely transfer of $ 5.5 billion worth of goods every day[2]. Many major ports are located in seismically active regions, and they are susceptibleto damage and disruption of service after a major earthquake. Resilience to full capacityof port facilities is necessary to prevent long-term losses which may result from tenantrelocation.

Marginal wharves are structures constructed on the shores of harbors that allow dockingof cargo ships for loading and unloading of goods in marine ports. Typical marginalwharves consist of piles driven into embankments of soil to the depth necessary toachieve the required compressive load capacity. The piles are then cast integrally withthe overlying concrete deck. Figure 1 provides schematics for a typical port plan view aswell as a typical marginal wharf elevation.

∗Corresponding author; e-mail: [email protected]†University of Washington, M.S. 2011§Associate Professor, Department of Civil and Environmental Engineering, University of Washington, Seattle‡Professor, Department of Civil and Environmental Engineering, University of Washington, Seattle

Copyright c© 2011 John Wiley & Sons, Ltd.

2 M.M. CHIARAMONTE ET AL.

Shipping Containers

Marshaling Yard

Shoreline

Container Crane

Marginal WharfCrane Rails

(a) Port plan view

Cra

ne

Leg

Cra

ne

Leg

or Rock Dike

Precast Driven Piles

Existing Sub−Surface Conditions

Cutoff Wall

Rock Fill & Rip−Rap

Back Fill

Crane Rail Concrete DeckCrane Rail Girder

(b) Typical wharf elevation

Figure 1. Marginal wharves in marine ports

Wharf facilities subjected to seismic demands must sustain forces resulting frominertial demands of the structure as well as kinematic demands from the embankmentdeformations. Lateral forces are resisted with battered piles or using plumb-pile systemswith moment resisting pile-to-wharf connections. Poor performance of some batter pilesystems have caused them to fall out of favor [9]. Plumb-pile systems are predominantlyused as the lateral force resisting system for most West coast U.S. ports. With plumb-pilesystems there are two primary locations of damage: at the pile-to-wharf connection andwithin the portion of the pile that is embedded in the soil. Damage at these locations caninitiate with small deformations (at deck displacements of ≈ 5% of the pile diameter).

Recent efforts in improving conventional pile-to-wharf connection design by Jellin [11]and Stringer [27] have resulted in a damage resistant pile-to-wharf connection design.The study presented in this paper sought to develop a keen understanding of traditionalmarginal wharf performance and of the impacts, on the overall system behavior, of theimproved damage resistant pile-to-wharf connection.

For this purpose finite element models for conventional and improved pile-to-wharfconnections were developed and validated against test data provided by Jellin [11]and Stringer [27]. Along with the connection models, piles’ sectional responses werevalidated with test data for several specimens from three research programs.

The validated piles and connections models were incorporated in a two dimensionalmodel of a typical section of a reference facility. To accurately simulate the seismicdemands on the structure the complete system was modeled. For a marginal wharf thisincludes a validated structural model as well as a soil continuum model with soil-structureinteraction springs. Although such approach is computationally demanding, Roth et al.[25, 24] have shown the need of a complete system model to properly capture demandson the structure from both inertia and soil kinematics. To further support the need of acomplete system model the work of Donahue et al. [6] investigated the applicability andaccuracy of various modeling techniques for marginal wharves and showed the need of afully coupled soil-structure model to properly capture the system’s response.

The complete finite element models were subjected to static pushover analyses as wellas dynamic time history analyses. A suite of thirteen recorded ground accelerations,scaled to three different hazard levels, was used for the dynamic time history analyses.

From static pushover as well as dynamic time history analyses, conclusions were drawnon the performance (structural damage and crane downtime) of marginal wharves duringseismic events and on the influence of the improved connection on the wharf response.Furthermore, through the use of probabilistic functions, estimates of port downtimeassociated with each hazard levels were evaluated and are herein summarized. It is worthnoting that analogous work has been carried out in [30] where fragility curves for wharfstructures are developed through the use of analytical results.

Copyright c© 2011 John Wiley & Sons, Ltd. Earthquake Engng Struct. Dyn. (2011)DOI: 10.1002/eqe

SEISMIC ANALYSIS OF CONVENTIONAL AND IMPROVED MARGINAL WHARVES 3

2. SIMULATED FACILITY

Following a review of fourteen major marginal wharf facilities on the U.S. west coast, amarginal wharf structure with typical properties was chosen as a representative facility tobe simulated . The chosen berth is situated in the San Francisco Bay. Figure 2 illustratesthe geometry of the chosen berth. The structure consists of a reinforced concreted decksupported by 61cm octagonal prestressed, precast concrete piles with moment-resistingembedded-dowel connections. A typical section, shaded in gray in Figure 2b, repeatsitself every 5.5m. Table I provides a summary of the pile properties at each pile row.A soil profile typical of the area was assumed. The chosen soil layering, along with itselevations, is provided in Figure 2c. The subsurface profile was chosen in order to be

2.1m

2.6m 4.6m 4 @ 5.5m 3.6m 2.6m

El. −32.3m

El. −39.3m

El. −15.2m

El. −13.7m

El. −12.2m

El. −10.7m

El. 4.6mE D C B A

0.9m 0.6m

4m

H G F

(a) Wharf Elevation View

A

B

C

E

F

G

H

D

2.4

m3

.6m

2.6

m Water Side2D − Planar Section

Land SideTyp

39

.4m

4 @

5.5

m4

.6m

4m

5.5m

(b) Wharf Plan View

Medium Dense Clayey Sand (Young Bay Deposit and S.A. Form)

El −6.1m

El 4.6m

El 0.0m

El −0.6m

El −4.0m

El −23.5m

El −61.0m

1.5

Stiff to Hard Clay (Old Bay Mud)

Dense to Very Dense Sand (San Antonio Formation)

Soft to Medium Stiff Clay (Young Bay Mud)

Submerged Rock Fill

Dry Rock Fill

(c) Subsurface Profile Elevations

Figure 2. Geometry and soil profile of simulated facility

consistent with the location of the reference facility.

3. FINITE ELEMENT MODEL

The typical section of the prototype facility was modeled in a two dimensional domainusing OpenSees [18] as the numerical platform. The analytical model is made up ofthree main components: the structure, the soil continuum and the discrete soil-structureinteraction (SSI) model.



Table I. Pile schedule

Dist. from top Spiral Spiral Bar No. of No. of PileRow Ranges (m) † Size Pitch (cm) Size Bars Strands Prestress (MPa)

A & B0-0.91 W11 3.8 No. 10 4 16 7.70.91-1.27 W11 7.6 No. 10 4 16 7.71.27-end W11 7.6 – 0 16 7.7

C & D0-0.91 W11 3.8 No. 10 4 14 6.80.91-1.27 W11 7.6 No. 10 4 14 6.81.27-end W11 7.6 – 0 14 6.8

E0-0.91 No.3 7.6 No. 10 4 14 6.80.91-1.27 W11 7.6 No. 10 4 14 6.81.27-end W11 7.6 – 0 14 6.8

F, G & H0-1.83 No.4 3.8 No. 10 8 16 7.71.83-2.29 No.4 6.3 No. 10 8 16 7.72.29- 7.60 W11 6.3 – 0 16 7.77.60-end W11 7.6 – 0 16 7.7

†- Ranges of pile lengths from pile tip

3.1. Structural Model

The finite element model of the marginal wharf structure consists of deck elements, pileelements and pile-to-wharf connections. The deck elements were assumed to remainelastic and modeled with elastic frame elements.

The piles were modeled using flexibility based, non-linear beam column elementsavailable in OpenSees. A total of eight pile rows with eleven piles support the wharfdeck. To address the reduced pile spacing at pile rows B, G and H, multiple elementsare overlaid. Specifically, at rows G and B for each pair of nodes two identical elementsconnect those nodes and at row H three identical elements are overlaid between nodes.

Within the pile elements each integration point is assigned a fiber cross sectionconsisting of unconfined and confined concrete, prestressing steel fibers as well as dowelbar fibers. The uniaxial response of the concrete was approximated by the Hognestadparabola to the peak strength (f ′c) and linearly softening post peak. Loading and reloadingfollows the response suggested by Sima et al. [26]. The mild steel constitutive relationfollows the Menegotto and Pinto [19] model superimposed with a Mansons and Hirshberg[16] damage model to capture fatigue induced fracture. The prestressing steel responseis approximated by the Menegotto-Pinto as well and an initial stress equivalent to itsprestress is assigned to it. Thus the cross-sectional response is capable of capturingconcrete failure as well as mild steel yielding and fracturing.

The sectional properties (i.e. content of mild steel, prestressing steel level ofconfinement) for each integration point depend on the location of the fiber section inthe pile element as defined in Table I. A total of thirteen fiber sections were developed tocapture the various cross sectional properties of the piles along their height and at theirrespective rows. Figure 3 illustrates a typical fiber section discretization along with itsdimensions.

To circumvent curvature localization issues associated with softening moment-curvature response of the force formulated frame elements the material constitutiveresponse was normalized to assure an objective response as described by Chiaramonte[3].

The response of the fiber section was validated through the use of moment-curvature comparisons with experimental data obtained from various test programs[8, 22, 12]. In Figure 4, two moment-curvature responses are compared to the respectiveexperimental response. Geometric non-linearities of the frame elements are addressed


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61

cm4

6 c

m

Fiber

Dowel

Strand

Fiber

Figure 3. Typical fiber section

0.0 0.1 0.2 0.3 0.4 0.5 0.60

100

200

300

400

Curvaure (rad/m)

Moment(kN

×m)

SimulatedExperimental

(a) Falconer - Unit 3

0.0 0.1 0.2 0.30

100

200

300

400

Curvaure (rad/m)

Moment(kN

×m)

SimulatedExperimental

(b) Joen - Unit 1

Figure 4. Selected moment curvature comparisons

with a corotationl transformation following the assumption of small deformations, largedisplacements.

Two connection types, referred to as conventional and controlled rocking (CR)connections, were investigated. The conventional dowel connection uses headed dowelbars to connect the wharf deck to the piles. In current practice, the dowel bars are groutedin galvanized steel ducts in the pile and cast integrally with the overlying deck. Figure 5ashows a diagram of the conventional dowel connection. These connections are susceptibleto early damage resulting from pile edge bearing. Once pile spalling takes place, theconnections loose considerable moment carrying capacity resulting in quick deterioration.

The improved CR connection was developed to mitigate the damage and deteriorationof the conventional connection [11, 27]. The CR connection incorporates the use of anannular bearing pad, a CERAMAR flexible expansion material around the perimeter ofthe pile and a 38cm long debonded section centered at the pile-deck interface. Two typesof annular bearing pads were used in the numerical simulations. One bearing pads was a1.9cm thick cotton duck (Sorbtex [29]) material. The second was a 1.3cm thick syntheticrubber (Fiberlast [29]) pad. Figure 5b shows a diagram of the improved connection detail.The improved CR connections reduce edge bearing stresses, which considerably delaydamage, and reduce deformation demands on the dowel bars, which results in delayedbar fracture.

The connection was modeled through the use of the validated pile element connectedto the deck through the use of zero length section elements. In OpenSees the zero lengthsection element represents a unit length element with a fiber section connecting two nodesof equal coordinates. The unit length property provides an element rotation equal to thefiber section curvature.



To capture the rotation at the pile-deck interface, the fiber section material properties inthe zero length element were defined by a stress-deformation response. The deformationresponse captured concrete crushing and dowel elongation from strain penetration andbar debond. The concrete bearing stress-deformation response was approximated by anHognestad parabola to peak strength and was assumed to be perfectly plastic post peakdue to the high level of confinement at the pile-deck interface. The slip response of thebonded rebar was modeled as suggested by Zhao and Sritharan [32], while the responseof the debonded dowel consists of a superposition in series of the model suggested byZhao and Sritharan [32] with a bilinear stress-elongation model of an elastic-perfectly-plastic dowel, where the elongation is computed as the strain (constant throughout thelength of the debond) multiplied by the debonded length. Furthermore the annular portionof the concrete fibers surrounding the core in the CR connection are replaced by padfibers with their respective stress-deformation response. The stress deformation responseis approximated by a trilinear model fitted to the convex stress strain function measuredthrough experimental tests. For more details on the constitutive laws used within theanalytical model the reader is referred to [3].

The connection response was validated against four specimens [11, 27]. Figure 5 showscomparisons between simulated hysteretic responses and their respective experimentalresults.

Rayleigh damping is introduced and is set to five percent of critical damping for thefirst and third mode of the wharf structure alone (i.e. disconnected from the soil).

3.2. Soil Continuum Model

A two dimensional plane strain soil continuum mesh was incorporated in the analyticalmodel to capture the wave dynamics and lateral slope movement from seismic excitations.The soil continuum model used a total stress analyses approach to capture the effects ofsoil dynamics. In this study the effects of liquefaction were purposely not considered inthe intent of reproducing the selected site existing conditions.

The soil continuum was modeled using a quadrilateral structured mesh. The depth ofthe soil continuum mesh was selected to accurately model the subsurface conditions tobedrock or to a substantially stiffer material than the overlying layers. Figure 6 illustratesthe geometry and discretization of the mesh. The distance from the slope to the meshboundaries was chosen to prevent boundary effects on the response of the embankmentat the location of the wharf. The depth of the soil mesh varied from 66m at the leftmostboundary to 44m at the rightmost boundary with a total width of approximately 389m.

The height of the quadrilateral elements was selected based upon the shear wave lengthof the softest material. To properly capture the shear wave propagation through the mesh,there must exist at least four elements for the shortest wave length to prevent relevanthigher frequencies from being filtered out. In order to estimate the shortest wave length,which controlled the required element height, from the frequency spectra of the groundmotions it was determined by observation that the maximum frequency content (fmax)in the ground motion input was 40 Hz. The maximum permissible height of the element(helt) was computed using Equation 1.

helt ≤λ

4=Tmin vmin

4(1)

where Tmin is f−1max and vmin is the shear wave velocity of the softest layer. A maximumelement height of 4 ft was determined.

Furthermore, the minimum time step for the dynamic time history analysis was takenas

dtmin ≤heltvmax

(2)



A − A

Deck

Grouted T−headed Bar

A − A

Pile

⇓(a) Conventional connection

A − A

Deck

A − AAnnular Bearing

Grouted T−headed Bar

Debonded Region

Pile

Padexpansion materialCERAMAR

⇓(b) CR connection

−10 −5 0 5 10−1.5

−1.0

−0.5

0.0

0.5

1.0

1.5

Rotation (% rad)

Moment(M

N×

m)

(c) Experimental conventional connection

−10 −5 0 5 10−1.5

−1.0

−0.5

0.0

0.5

1.0

1.5

Rotation (% rad)

Moment(M

N×

m)

(d) Experimental CR connection

−10 −5 0 5 10−1.5

−1.0

−0.5

0.0

0.5

1.0

1.5

Rotation (% rad)

Moment(M

N×

m)

(e) Simulated conventional connection

−10 −5 0 5 10−1.5

−1.0

−0.5

0.0

0.5

1.0

1.5

Rotation (% rad)

Moment(M

N×

m)

(f) Simulated CR connection

Figure 5. Selected moment rotation comparisons of experimental results with simulations

389m

66m

44m

Figure 6. Finite element mesh



to prevent the shear waves from skipping elements during the analysis, where vmax isthe shear wave velocity of the stiffest material. A minimum time step of 0.002 sec. wasconservatively chosen.

With only a portion of the subsurface being modeled, special care was taken inmodeling the boundary between the soil continuum and the underlying finitely rigid halfspace. In fact, if the boundary is not properly handled, shear wave energy may remaintrapped vertically in the analytical model. This would mimic the contact between themodeled soil and an infinitely rigid layer (fixed boundary).

To account for the finite rigidity of the underlying layer the ground motion was input asa force, FI , coupled with a dashpot damper, of coefficient c, as suggested by Joyner andChen [13] and Lysmer and Kuhlemeyer [15], where

FI = ρEvs2u̇IA (3)

c = ρEvsA (4)

in which ρE and vs are the mass density and shear wave velocity of the underlying elastichalf space respectively. A is the cross sectional area of the mesh at the boundary and u̇Iis the particle velocity of the incident wave. This procedure prevented wave energy frombeing trapped within the analytical model and provided an accurate prediction of the wavepropagation through the soil continuum. Note that if the control motion is a rock outcroprecord, the recorded velocity is equivalent to 2u̇I .

The lateral boundaries consisted of two soil columns to the left and right of thesoil mesh. The purpose of the soil columns was to capture the horizontal propagationof seismic waves. The quadrilateral elements’ thickness, for the soil columns wasconsiderably larger than the one used for the interior mesh to ensure that the free fieldresponse was not affected by the interior soil response. Furthermore, to ensure that thefree-field columns captured solely the shear wave propagation, the nodes of the sameelevation were restrained to one another in both degrees of freedom allowing solely pureshear deformation modes. Figure 7 illustrates the schematic of the continuum soil model.

EqualDOF

Free field column

EqualDOF

Free field column

c = ρEvsAbase

F = ρEvsAbase2u̇I

Figure 7. Schematic of continuum soil model

Material nonlinearity for the cohesive and non-cohesive soils was capturedwith the help of two material models, the PressureDependentMultiYield andPressuIndependentMultiYield constitutive models [7, 31], which are available in


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Table II. Soil constitutive properties

Soil Type φ ρs G K γmax p c(ton/m3) (MPa) (MPa) (kPa) (kPa)

Dry Rock Fill 45 ◦ 2.0 210.0 340.0 0.1 82.7 –Submerged Rock Fill 45 ◦ 2.4 210.3 340.0 0.1 82.7 –Young Bay Mud – 1.9 33.6 165.5 0.1 – 27.6Young Bay Deposit & SanAntonio Formation

– 2.2 85.8 290.0 0.1 – 48.3

San Antonio Formation 42 ◦ 2.4 180.0 310.0 0.1 82.7 –Old Bay Mud – 2.0 310.0 606.7 0.1 – 206.8

φ - Friction Angleρs - DensityG - Shear ModulusK - Bulk Modulusγmax - Peak Shear Strainp - Reference Pressuren - Pressure Dependent Coefficientc - Cohesion at Zero Confinement

OpenSees. Table II provides a summary of the input parameters used for the constitutivemodels. The values of the parameters were based on recommendations available in theOpenSees manual [18] for the selected site soil conditions.

In these models the cohesive yield criterion is defined by a one invariant (J2) von Misestype surface and the constitutive material nonlinear hardening response is defined by thetranslation of n nested surfaces, with respective tangent properties, inside the boundingfailure surface. For the cohesive soils, plasticity is of the associative type.

For non-cohesive soils the model is defined by a two invariant (I1,J2) Drucker-Pragertype yield criterion. Similarly to the cohesive soils, the non-linear response inside thebounding surface is defined by n nested surfaces. In these case, plasticity is of thenon-associative type and accounts for volumetric contraction and dilation resulting fromshearing. To properly capture the pressure dependent inelastic properties, an elasticgravity analysis was conducted, after which the material state was updated to inelastic.The elastic gravity analysis developed the overburden pressures necessary for the pressuredependent failure surface to become of finite dimensions in deviatoric space. The gravityanalysis was taken as instantaneous and no considerable deformations were observed inthe updating the material from elastic to inelastic state.

No artificial damping was introduce for the soil model as all of its dissipative propertiesare believed to originate from material non-linearities.

3.3. Soil-Structure Interaction Model

Soil structure interaction was modeled using non-linear uniaxial springs as illustratedin Figure 8. The distributed uniaxial non-linear springs represent a discretization of thecontinuous pile-soil interaction. The springs connect a two-dimensional solid element to athree-dimensional beam element and account for the three-dimensional pile-soil interfaceinteraction response. Properties of each uniaxial spring are intended to represent theresponse of the springs tributary width. The springs are of three types: p-y, t-z and Q-z springs. P-y springs simulate the lateral resistance of the soil surrounding the pile alongits height, t-z springs simulate the vertical resistance resulting from skin friction alongthe height of the pile and Q-z springs provide the bearing resistance of the soil at the pileend. The p-y springs followed the response suggested in API [1] and in Matlock [17] forsands and clay respectively. The t-z springs followed the response suggested in Mosher



P−y spring

Q−z spring

Soil Mesh

Equal D.O.F. w/ Soil Meshor Fixed

Pile Node

Pile Element

N

V

M

N

V

M

T−z spring

Figure 8. Soil Springs

[20] and in Reese and O’Neil [23] for sands and clay respectively. Furthermore the non-linear behavior of the Q-z springs followed the response suggested in Vijayvergiya [28]and Reese and O’Neil [23] for sands and clays respectively.

The p-y and t-z springs were spaced 30cm apart along the height of the piles. The Q-zsprings were placed at the base of the pile. Soil springs were modeled using a zero lengthelement with a master node shared with the pile element and a slave node connected tothe 2D soil mesh (during dynamic analyses) or fixed in space (during pushover analyses).The properties of the discrete soil springs were computed using the existing subsurfaceprofile at the location of the springs.

Similarly to the soil model, no artificial damping was introduce for the soil-structureinteraction model as all of its dissipative properties are believed to originate from materialnon-linearities

3.4. Model Constraints

The soil continuum was intended to enable the accurate simulation of the dynamicresponse of soils. For the static pushover analysis the soil continuum was not needed,as its sole purpose is to capture the seismic wave propagation, therefore the springs’ slavenodes were restrained in all of their degrees of freedom.

For the dynamic analyses, when a soil continuum was generated along with thestructural and the SSI model, the slave nodes of the soil springs were constrained tothe soil continuum mesh. Generating a soil mesh where for each spring node existsa soil mesh node may result in an inefficient and expensive meshing. To reduce thecomputational demand of the numerical model, it was assumed reasonable to constrainthe soil springs to the closest soil mesh node. This allowed the generation of the soilmesh to be fully independent of the SSI spring nodes and for space to exist between theconstrained springs’ slave node to its counterpart in the soil mesh. Using this approach,more than one spring node may be constrained to a single mesh node, if the latter wasthe closest available. The mean distance in the analytical model between the two nodesfor all the springs was approximately 31cm with a maximum distance of 1.3m. Althoughthe constraints may not be exact, it was assumed that the effects on the global responseof the wharf facility would be negligible, while the overall savings, in terms of degrees offreedom, substantial.



Table III. Ground motions summary

NGA # Name Moment PGA VS30 Joyner-BooreMagnitude (g) (m/s) Distance (km)

NGA0072 San Fernando 6.61 0.19 821 19.5NGA0080 San Fernando 6.61 0.20 969 21.5NGA0146 Coyote Lake 5.74 0.13 1428 10.7NGA0663 Whittier Narrows 5.99 0.18 821 14.5NGA0680 Whittier Narrows 5.99 0.11 969 6.8NGA0765 Loma Prieta 6.93 0.47 1428 8.8NGA0957 Northridge 6.69 0.16 822 15.8NGA0994 Northridge 6.69 0.29 1015 21.2NGA1011 Northridge 6.69 0.17 1222 15.1NGA1041 Northridge 6.69 0.23 821 35.3NGA1091 Northridge 6.69 0.15 996 23.1NGA1645 Sierra Madre 5.61 0.28 822 2.6NGA1649 Sierra Madre 5.61 0.12 996 37.6

To avoid gravity induced settlement deformation demands on the piles and wharf, thesprings’ slave nodes were restrained to their closest nodes in the soil mesh only after thesoil gravity analysis was conducted.

4. GROUND MOTIONS

The ground motions were selected from the Next-Generation Attenuation of GroundMotions (NGA) database [4]. The following criteria were used for the selection:

� Minimum moment magnitude of 5.5� Closest distance to rupture 0-50 km� Minimum vs30 of 800 m/s� Strike-slip, reverse and reverse oblique mechanism� Located in coastal California

These ground motion parameters are representative of the port site. For each record,only the horizontal motion with the larger peak ground acceleration was used. Table IIIsummarizes the motions characteristics and Figure 9 shows the response spectra for thethirteen ground motions for a critical damping ratio of 5%. The ground motion componentchosen for the analysis was the one with the largest peak ground acceleration.

Three target response spectra were built for a 2% in 50 year event, a 5 % in 50 yearevent and a 10% in 50 year event. The values of Ss and S1 for each hazard level wereobtained from the 2008 U.S. Geological National Seismic Hazard Maps [21].

The same suite of ground motions were scaled so that the mean of the response spectraover the range 0.2T1 to 1.5T1 , with T1 being the period of the first mode of oscillationof the structure, was greater or equal to the mean of the target response spectra overthe specified range. The value of T1 was computed to be 0.173s for the conventionalconnection detail and 0.184s for the CR connection detail.

5. STRUCTURAL & GEOTECHNICAL PERFORMANCE

Three analytical models were subjected to static pushover loads and dynamic earthquakeexcitations. The analytical models differed only in connection types. Gravity loadsconsisted of one and a half the dead load of the structure as recommended in [5].The three



10−2

10−1

100

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T (s)

Sa(g)

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10−2

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Target Response SpectraMean of Scaled Response Spectra

0.2 Tn ≤ T ≤ 1.5 Tn

Tn

(b) Mean of scaled vs. target response spectra (2% in 50 yrs.p.e.)

Figure 9. Ground motions response spectra - ξ = 5%

pile-to-wharf connections consisted of a conventional connection, a controlled rockingconnection with a 1.9cm cotton duck pad (Sorbtex connection) and a controlled rockingconnection with a 1.3cm thick synthetic rubber pad (Fiberlast connection).

5.1. Static Pushover Analyses

A set of static pushover analyses were conducted to determine the system displacementscorresponding to several damage states, to evaluate the effects of soil structure interactionand to determine a plausible location of the inflection point. The analyses were conductedby applying a lateral load to the deck to achieve a target deck displacement. The ultimatetarget displacement was taken as 25.4cm.

The results of the pushover analysis show that connection spalling is considerablydelayed with the use of the improved connection, while pile body damage in thesoil occurs at fixed deck displacements, regardless of connection type. While with theconventional connection spalling initiates at a wharf displacement of approximately5cm, with the CR connections spalling is delayed to 18cm and 20cm displacementsfor the Sorbtex and Fiberlast pads respectively. Furthermore pile body damage occursat approximately 5cm displacements for all connections. Figure 10 illustrates the threepushover curves. Displacements corresponding to specific damage levels are indicated inthe Figure.

The results of pushover analyses provided means to estimate the location of theinflection point for the port investigated. Figure 11 shows the moment distributions alongthe height of the pile at different port displacement throughout the pushover. While somevariability of the location of the inflection points existed, it is reasonable to assume thatthe inflection point was approximately 1.2m to 1.5m (two to two and a half pile diameters)away from the connection for the conventional detail while it was shifted to 0.9 to 1.4m(one and a half to two and a half pile diameters) from the connection for the CR detail.

The inflection point for the CR connection was closer to the deck than its conventionalcounterpart. This was due to additional flexibility of the CR connection, as seen in Figure5, a result of the debonded dowels and the bearing pad configuration.

As previously described, the soil-structure interaction was modeled using of uniaxialp-y, t-z and q-z springs. The lateral resistance p-y springs simulated the soil reactionagainst the pile. Figure 12 illustrates the spring axial forces and axial deformations alongthe height of a pile at row H. The multiple lines correspond to the incrementing deckdisplacement. Only the top seven lateral springs (top two meter of the subsurface) of row



0 5 10 15 20 250

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initiates ~ 6cm

spalling initiates~ 17cm

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spalling initiates

~ 20cmspalling initiatesFiberlast conn.

~ 5cm

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Figure 10. Port static pushover curves

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~1.35m~1.15m

Figure 11. Moment distributions in pile row H

H fully yielded during the pushover analysis. The above suggests a possible depth to fixityof approximately 3 to 4 pile diameters or 2.4m.

−1 −0.5 0 0.5 1 1.5−40

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Deformation (cm)

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Yielded Springs

(b) P-y springs axial deformations

Figure 12. Pushover p-y springs axial response along height of a pile at row H

5.2. Dynamic Time History Analyses

The wharf model with the three varying connections was subjected to a suite of thirteenground motions scaled to the 2500, 1000 and 500 year events. Figure 13 shows that eventhe demands from the larger event (2 % P.E. in 50 years) were not sufficient to causespalling for the three connection types and for the selected soil conditions. The wharfdeformations did not change substantially amongst connection types. The above suggeststhat the CR connection does not drastically affects the response of the marginal wharfalthough it is capable of sustaining considerably larger rotation demands with no damage.

Considerable deformation was evident in the soft clay (Young Bay Mud) layer asshown in Figure 14. The localization of deformation in the soil mesh is largely dependenton the element height. For the element size used within the model, the localization ofdeformation in the clay layer did not lead to severe curvature demands on the pilesexceeding the ones associated with spalling initiation.

While the localization of deformation in the soft clay did not lead to considerabledamage, its large inelastic deformations may have contributed in dampening the surfacemotion reducing the inertial demands on the structure.



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rad)

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ConventionalCR -SorbtexCR- Fiberlast

(c) 2% P.E. in 50 yrs events

Figure 13. Maximum connection rotation demands

6. NON-STRUCTURAL PERFORMANCE

The recorded acceleration time histories measured at the deck surface (from the numericalanalyses) were used in combination with probabilistic tools to estimate crane damage andport downtime.

To evaluate the response of cranes and the downtime associated with crane damage,acceleration response spectra were generated from the accelerations recorded at the deckwith a critical damping ratio of ξ = 1.5% chosen following the observations of Jacobs[10] and Kosbab [14]. The computed spectral accelerations were compared with fragilityfunctions (with spectral acceleration as the engineering demand parameter) to estimatethe probability of occurrence of a certain crane damage state and the port loss, in termsof time, associated with it.

Three crane types, representative of currently operating ones, were used for theprediction of down time. The crane types consisted of the jumbo J100, the LD100 (aprecursor of the modern jumbo cranes), and the LD50 (representative of older, smallcranes).

For each ground motion, connection type and hazard level the spectral accelerationresponses were computed and the corresponding spectral acceleration to the naturalperiods of each crane type were extrapolated. .The extrapolated spectral accelerations



y

x

x−displacements (cm)Young bay Mud

−9

9

0

Figure 14. Young Bay Mud shear band of permanent deformation

were used in combination with fragility functions developed by Kosbab [14] to estimatecrane damage and port downtime. For the 500 year event the port downtime was estimatedto be as long as 8 months depending on the ground motion while for the 1000 yearsevent the downtime was estimated to be over 9 months. As expected the largest loss wasdetermined to be associated with the 2500 years event where the expected downtimeassociated is considerable for most motions and crane types, with maximum downtimesestimated of over 10 months as shown in Figure 15.

The drastic change in response between the LD100,J100 and LD50 cranes can beattributed to the older nature of LD50 cranes and to the considerable difference of thefrailty functions used. The interested reader is referred to [14]. Furthermore it can benoted that for some ground motions the crane downtime associated with the conventionalis less than the one associated with the CR connection. The above is attributed tothe increased flexibility introduced by the CR connection which thus leads to largeracceleration demands on the cranes.

7. SUMMARY AND CONCLUSIONS

To evaluate the seismic performance and the effects on wharf response of the recentlydeveloped damage resistant (or controlled rocking) pile-to-wharf connection, a twodimensional finite element model of a reference structure was developed and analyzedconsidering static and dynamic loading. The numerical simulations modeled a completesystem which included a detailed soil and soil-structure interaction model along withthe structural model. Such approach was deemed necessary under the considerationthat neglecting the soil dynamic response would preclude considerable demands on thestructure from soil dynamics and kinematics.

A reference facility with properties typical amongst fourteen operating wharfstructures, was identified. Finite element models for conventional and controlled rockingconnections, as well as for pile elements, were developed and validated against test data.The models of the structural components were then implemented in the analytical modelof the marginal wharf.



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LD50ConventionalCR - SorbtexCR -Fiberlast

(c) LD50 cranes

Figure 15. Port loss in terms of time from 2% P.E. in 50 yrs event

Along with the structural model, a two dimensional, two degrees of freedom soilcontinuum mesh was included to capture the free field wave propagation response andlateral spreading, which, combined with the uniaxial soil structure interaction springs,provided a reasonable estimate of the soil contribution to the structural demands.

From the results of static pushover analyses, as well as thirteen time history analysesfor three hazard levels (2 %, 5%, 10% probability of exceedance in 50 years), conclusionswere drawn on the seismic performance of marginal wharves.

The static pushover analyses highlighted the superior performance of the recentlydeveloped controlled rocking connection over the conventional detail, as damage isconsiderably delayed if not fully mitigated.

During dynamic time history analyses the demands from inertial forces alone wasinsufficient to cause damage in the facility. The demands from soil deformations werenot substantial.

While little damage was observed in the marginal wharf structure throughout thedynamic time history analyses, with the use of probabilistic tools it was estimated thatthe loss associated with crane damage and downtime was substantial for most motionsand hazard levels.

The above work could be further extended to have a broader impact through a series ofparametric studies on the subsurface influence on the overall response of the structure aswell as the cranes.



8. ACKNOWLEDGMENTS

This work was supported by the National Science Foundation, under the award numbersCMS-0530478 and CMS-0402490 through the Georgia Institute of Technology. Anyopinions, finding, conclusions or recommendations expressed in this material are thoseof authors and do not necessarily reflect those of the National Science Foundation.

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Seismic analyses of conventional and improved marginal wharves

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